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Analysis of metallurgical defects in enamel steel castings

  • Qixuan Rui EMAIL logo , Zijian Huang , Yingjiang Li and Xiaoguang Hu
Published/Copyright: March 19, 2024
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High Temperature Materials and Processes
From the journal High Temperature Materials and Processes Volume 43 Issue 1

Abstract

The relevant experiments and studies were conducted to address the metallurgical defects such as subcutaneous bubbles and slag inclusions occurring during the enamel steel continuous casting process. According to the high-temperature experimental results, calculations were made to determine the changes in viscosity and tension due to the steel reaction with the slag. Based on the experimental findings, there was a notable variation in the content of SiO2 and Al2O3 in the slag before and after the reaction. The concentration of the element Al in the steel melt significantly decreased, whereas the concentration of Ti showed a minimal change. These observations indicate that the reduction of SiO2 at the slag–steel interface is predominantly attributed to the role of Al. The calculation results showed that at 300 s of reaction time, the viscosity rapidly increased from 0.108 to 0.133 Pa·s and then slowly increased to 0.155 Pa·s; The interfacial tension rapidly decreased from its initial value of 1,380 mJ·m−2 to a minimum of 1,320 mJ·m−2, and it then slowly increased to an equilibrium state. Therefore, the main cause of the occurrence of porosity defects in the enamel steel continuous casting process is the reduction of interfacial tension between slag, steel, and gas caused by the reaction between the slag and steel, as well as the foaming of the slag.

Keywords: surface defects; interface reaction; interfacial tension

1 Introduction

Cold-rolled enamel steel is a type of ultra-low carbon steel containing Ti, with a high content of elements such as Al, Ti, and N. In a certain steel plant, serious quality defects such as subsurface gas bubbles and inclusion were observed in the continuous casting process. Regarding the formation of subsurface gas bubbles and inclusion defects, the current main view is that argon gas injected through a stopper rod or an immersed nozzle will flow with the molten steel into the mold. The bubbles will rise to the surface of the slag in the mold, and when bubbles cannot escape from the top surface of the mold in time, they will be trapped by the solidified shell of the billet, resulting in subsurface gas bubble defects [1]. After high-speed injection at the nozzle outlet, argon gas inside the mold forms bubbles. When the gas flow rate is high, the argon gas bubbles impact the slag–steel interface, causing liquid surface fluctuations and resulting in slag entrapment, which leads to inclusion defects [2].

Relevant research shows that to obtain a lower melting point, the mold flux usually contains a higher content of reducible oxides such as silicon dioxide and iron oxide. When the [Al] content in the molten steel is high, it reacts with these oxides in the slag [3]. According to relevant literature, when the [Ti] and [N] contents in the molten steel are high, TiN inclusions are formed in the steel at the casting temperature. When these TiN inclusions come into contact with the mold flux, a reaction occurs, resulting in the formation of aluminum titanium oxide and nitrogen gas, ultimately leading to the “fish-eye” phenomenon [4].

This study focuses on the flux–steel reaction between enameled steel and CaO–Al2O3–SiO2–Na2O–CaF2–Fe2O3 slag system in a continuous casting process. The change of chemical composition of steel and slag after reaction with titanium-containing steel was analyzed by high-temperature experiment. The phase composition and viscosity of slag after the flux–steel reaction were calculated by FactSage 7.2 thermodynamic analysis software. The change of interfacial tension between flux and steel after flux–steel reaction is calculated by a theoretical model formula. This study is of great significance for improving the surface quality of enameled steel continuous casting billet.

2 On-site detection of metallurgical defects in cast billets

2.1 Electron microscopic examination of billet samples

Currently, the occurrence of gas porosity defects (including “bubbles,” “subsurface bubbles,” and “pinholes”) in the cold-rolled enamel steel slab produced by a certain steel plant’s continuous casting machine is quite frequent, significantly impacting the production of the slab production line. The surface gas porosity defects on the cold-rolled slab mainly appear as elongated, fish-eye-like pits, as shown in Figure 1. To determine the cause of the formation of gas porosity at the edges of the enamel steel billets, steel samples were specifically taken from the edges of the ingots where this phenomenon was observed on-site. The steel samples were ground and polished, and then, the pores and their surroundings were examined using an electron microscope. The results of the examination are shown in Figure 2.

Figure 1 Defects in enamel steel castings.
Figure 1

Defects in enamel steel castings.

Figure 2 Scanning electron microscopy-Energy dispersive X-ray spectroscopy (SEM-EDS) analysis of surface defects in cast billet samples: (a1, a2) SEM-EDS at larger pore point 1 and (b1, b2) SEM-EDS at small gas pore point 1.
Figure 2

Scanning electron microscopy-Energy dispersive X-ray spectroscopy (SEM-EDS) analysis of surface defects in cast billet samples: (a1, a2) SEM-EDS at larger pore point 1 and (b1, b2) SEM-EDS at small gas pore point 1.

2.2 Analysis of detection results

According to the detection results, examination of larger pores in the electron microscope shows that the pores and their surroundings are mainly composed of elements such as Al, Si, and Fe, as indicated by point scanning in Figure 2(a2). Examination of smaller pores in the electron microscope indicates that the composition around the small pores is mainly composed of elements such as Ti, N, and Fe, as indicated by point scanning in Figure 2(b2).

Based on the detection results, it is believed that therefore, it can be inferred that the larger pores are formed by argon bubbles introduced into the mold through tundish argon blowing. Since bubbles can adsorb inclusions, the region surrounding the larger pores mainly consists of various elements, primarily Al. This region represents the residual porosity that has adsorbed inclusions. On the other hand, the smaller pores are surrounded by elements such as Ti, N, and Si, indicating a possible reaction [5]:

(1) TiN(s) + SiO 2 (s) = SiTi(s) + [ Si ] + 1 / 2 N 2 (g) Δ G ( 1 ) 0 = 206 , 094 − 120.93 T ( J ⋅ mol − 1 ) .

Furthermore, according to the detection results, it is observed that the pores are also surrounded by a certain amount of Na element. Na is a typical component of the mold flux, indicating that SiO2 in the mold flux is reduced by [Ti], and this leads to alterations in both the composition and characteristics of the slag. This can also lead to the involvement of the slag [6].

Based on scanning electron microscopy and energy spectrum analysis, the following conjectures can be made:

  1. The larger pores are formed by argon bubbles introduced into the mold through tundish argon blowing.

  2. The smaller pores within the mold flux can be attributed to the reaction between TiN precipitates present in the steel and SiO2, releasing nitrogen gas and being carried into the solidified steel.

Therefore, through the analysis of surface defects on ingots in the enamel steel production site, it was found that the subcutaneous porosity defects in this type of steel are mainly caused by large inclusions in the steel, namely, “bubbles + Al2O3 + TiN”-type inclusions and inclusions formed by the entrapment of mold flux. It can be seen that the formation mechanism of these porosity defects is closely related to the reaction with steel and mold flux. To further clarify the impact of the reaction between steel and mold flux on the surface defects of enamel steel, how the steel and mold flux interface reaction affects surface defects will be studied in depth below. This will continue to be studied through high-temperature simulation experiments and thermodynamic analysis.

3 High-temperature experiments and experimental methods

3.1 Experimental materials

This experimental research focuses on cold-rolled enameled steel (Grade MBTTC20001). To comprehensively investigate the flux–steel reactions and their effects, the experiment utilized mold flux and steel with the same composition as those used in the scene. The experimental mold flux employed chemically pure reagents of identical composition. Table 1 provides the element content of experimental steel, while Table 2 presents the chemical compositions of the slag.

Table 1

Chemical composition of the steel used in the present experiment

Component C Si Mn P S Als Ti N
Content (wt%) ≤0.004 ≤0.03 0.10 ≤0.012 0.03 0.02–0.04 0.10–0.12 0.009
Table 2

Chemical compositions of the slag used in the present experiment

Component CaO SiO2 Al2O3 Na2O CaF2 Fe2O3
Content (wt%) 40 35 5 10.5 8 1.5

3.2 Experimental device

The experiments on slag–steel interface reactions were carried out using a vertical induction resistance furnace configuration. As illustrated in Figure 3, the apparatus mainly consists of an induction coil device, refractory materials, a bottom thermocouple, and a corundum reaction tube. Each experimental batch involved 1,000 g of steel, prepared using industrial-grade pure iron as the raw material. The mold flux quantity for each batch was 50 g, and the raw materials used were chemically pure reagents following the composition shown in Table 2. The chemical pure reagents, including CaO, SiO2, Na2O, Al2O3, Fe2O3, and CaF2, were uniformly mixed in a sealed bag. During alloy smelting, 50 g of slag was loaded into an alumina crucible, which was enclosed within a graphite crucible for protection, and placed in the constant-temperature zone of the corundum tube. Before the start of the experiment, high-purity argon gas was introduced from the top of the furnace to prevent oxidation of the steel. Efforts were made to remove oxygen from the tube and maintain continuous argon gas protection throughout the process.

Figure 3 Schematic diagram of induction resistance furnace device.
Figure 3

Schematic diagram of induction resistance furnace device.

3.3 Experimentation

Before conducting the experiment, a precise mixture of industrial-grade pure iron and titanium nitride powder was prepared and deposited into a high-purity alumina crucible. The crucible was surrounded by a graphite crucible and placed in the constant-temperature zone of the corundum reaction tube. To safeguard the steel from oxidation, continuous argon gas protection was maintained throughout the experiment, with a flow rate of 0.3 L·min−1 of high-purity Ar introduced for at least 30 min to purge any oxygen present during the reaction process. After initiating the power supply for temperature ramping, the flow rate was adjusted to 2 L·min−1, and the temperature was increased at a rate of 4℃/min until reaching 1,550℃, followed by a 20 min holding period to ensure complete melting of the experimental steel sample. The prepared pre-melted mold flux material was rapidly added to the surface of the steel in the alumina crucible through a quartz tube, and the reaction time was recorded from the onset of the reaction. Subsequently, the melting was carried out in the induction furnace. The smelting temperature and holding time were set at 1,823 K (1,550°C) and 30 min, respectively, to ensure uniformity in chemical composition and sufficient reaction at the mold flux and steel. When the mold flux and steel samples in the crucible at the center of the furnace begin to contact and react, through the observation window on the surface of the furnace, an optical camera is used to observe the dynamic changes of the flux–steel interface on the surface of the sample captured at different reaction time intervals. Then, combined with computer microscope software, measure and evaluate the surface contact angle between molten steel and slag at different times. Take out the steel and slag samples from the furnace at 5, 10, 15, 20, and 30 min, respectively, and rapidly quench them in water. The slag sample is taken at the flux–steel interface, and the steel sample is taken at the steel–steel interface. Furthermore, grind the steel and slag samples into powder form and mark them with corresponding numbers. The chemical composition of slag samples was determined by X-ray fluorescence spectroscopy analysis, and the elemental content of Al, Ti, and Si in steel samples was determined by inductively coupled plasma mass spectrometry analysis. The bulk steel and mold flux samples were taken at the steel interface and slag interface, respectively, and the interface phenomena at the flux–steel interface were analyzed by SEM-EDS, and the phase composition of the mold flux samples after reaction was analyzed by XRD.

4 Experimental results and discussion

4.1 The basic law of changes in the components of mold flux and steel caused by the interface reaction between mold flux and steel

To elucidate the interaction between the slag and enameled steel, the compositional changes of the slag were analyzed at different reaction times. By analyzing the reaction mechanism between slag and steel from a thermodynamic perspective, samples were collected and analyzed at various reaction times under the condition of 1,550°C. Record the test results of the slag and steel samples as A0–A5 and B0–B5, respectively. The compositional changes of the mold flux are presented in Table 3 and Figure 4, while the variations in [Al], [Si], and [Ti] components in the steel melt are shown in Table 4 and Figure 5.

Table 3

Chemical composition changes of slag at different reaction times (mass%)

Number t (min) CaO Al2O3 Fe2O3 Na2O CaF2 SiO2 TiO2
A0 0 40 5 1.5 10.5 8 35 0
A1 5 39.09 7.41 1.13 10.42 7.95 32.69 1.23
A2 10 38.57 8.57 0.88 10.38 7.94 29.97 1.68
A3 15 37.92 9.02 0.65 10.23 7.95 27.88 1.79
A4 20 37.02 9.65 0.51 10.15 7.93 25.78 1.94
A5 30 36.88 9.95 0.42 10.08 7.93 25.10 1.96
Figure 4 Changes in composition of slag with reaction time of slag and steel: (a) changes in CaO, Al2O3, and SiO2 over time and (b) changes in TiO2 and Fe2O3 over time.
Figure 4

Changes in composition of slag with reaction time of slag and steel: (a) changes in CaO, Al2O3, and SiO2 over time and (b) changes in TiO2 and Fe2O3 over time.

Table 4

Changes in [Al], [Si], and [Ti] elements in steel at different reaction times (mass%)

Number t (min) [Al] [Si] [Ti]
B0 0 0.78 0.4 0.52
B1 5 0.35 0.65 0.45
B2 10 0.2 0.98 0.36
B3 15 0.1 1.3 0.13
B4 20 0.08 1.35 0.1
B5 30 0.05 1.4 0.08
Figure 5 Changes in [Al], [Si], and [Ti] elements in steel.
Figure 5

Changes in [Al], [Si], and [Ti] elements in steel.

Figure 4(a) shows that with the extension of reaction time between liquid steel and mold flux, the content of (SiO2) in slag steadily decreases from 35 to 25.10%, while the content of (Al2O3) increases from 5 to 9.95%. Correspondingly, as shown in Figure 5, the content of [Al] in molten steel is significantly reduced due to the reaction between molten steel and mold flux, from 0.78 to 0.05%, while the variation in [Ti] content changes is relatively minor. It indicates that [Al] reduction (SiO2) is the main reaction at the flux–steel interface, while [Ti] has little effect on the accumulation of (Al2O3) in slag. Therefore, the reduction in [Al] in the liquid steel, the decrease in (SiO2) content in the slag, and the increase in (Al2O3) content can mainly be attributed to chemical reactions (2) taking place at the interface between the steel and slag. In the initial stage of the reaction (<5 min), the slag–steel reaction is intense, leading to rapid changes in the chemical composition of the steel melt and the slag. However, after 5 min of reaction, the [Al] content in the steel melt and the SiO2 content in the slag gradually stabilized. This suggests that the oxidation reaction of [Al] in the steel melt and the reduction reaction of SiO2 in the mold flux may predominantly occur within the first 5 min of the reaction.

(2) 3 ( SiO 2 ) + 4 [ Al ] = 2 ( Al 2 O 3 ) + 3 [ Si ] Δ G ( 2 ) 0 = − 746 , 400 + 114.48 T ( J ⋅ mol − 1 )

As shown in Figure 4(b), during the initial stage of the reaction (0–5 min), the TiO2 content in the flux rapidly increased from 0 to 1.23%. This increase is primarily attributed to the reaction between SiO2 in the slag and Ti in the liquid steel at the flux–steel interface (3). Therefore, this reaction leads to a decrease in Ti content in the steel and an increase in TiO2 content in the slag. In the intermediate stage of the reaction (5–15 min), there continues to be a gradual increase in TiO2 content in the slag, albeit at a slower rate. Chen [7] conducted thermodynamic analysis, which revealed that TiN inclusions present in titanium-containing steel react with SiO2 and Fe2O3 present within the slag (reactions (4) and (5)). As a result, this leads to a reduction of Fe2O3 content within the slag while increasing its TiO2 content. Additionally, Al2O3 within the steel may combine with TiO2 with the slag, resulting in the formation of titanium–aluminum complex inclusions at the flux–steel interface. This is also the reason why the change rate of TiO2 content slows down. Relevant study [8] has found the presence of binary phase (TiO2·Al2O3) in titanium-containing low-carbon steel by experimental study. It also confirms this view. In the later stage of the reaction (15–30 min), the (TiO2) content in the slag gradually stabilizes but still exhibits a slight increasing trend. This can be attributed to the generation of a significant number of nitrogen bubbles at the interface due to the chemical reaction [9], and titanium oxide inclusions float to the slag under the action of the buoyancy of the nitrogen bubble, resulting in a slight increase in (TiO2) content.

(3) [ Ti ] + ( SiO 2 ) = ( TiO 2 ) + [ Si ] Δ G ( 3 ) 0 = − 101 , 050 + 7.3 T ( J ⋅ mol − 1 ) ,

(4) TiN + 2 ( Fe 2 O 3 ) = ( TiO 2 ) + 2 Fe ( 1 ) Δ G ( 4 ) 0 = − 403 , 780 + 115.19 T ( J ⋅ mol − 1 ) ,

(5) SiO 2 + TiN = Si ( s ) + TiO 2 + 1 / 2 N 2 ( g ) Δ G ( 5 ) 0 = 206 , 094 − 120.93 T ( J ⋅ mol − 1 ) .

4.2 Effect of flux–steel reaction on the viscosity of mold flux

In order to study the influence of chemical composition and phase change of mold flux caused by the flux–steel reaction on its viscosity, the CaO-SiO2-Al2O3-TiO2-Na2O-Fe2O3 (1.23% TiO2, 10.42% Na2O, 1.13% Fe2O3) phase diagram was qualitatively calculated by FactSage7.2 thermodynamic software. As shown in Figure 6, it can be found that the slag components rapidly approach the region with higher melting points. As can be seen from Figure 4 2(a), the contents of (SiO2), (CaO), and (Al2O3) in the slag all change with the passage of reaction time, so the basicity of the protective slag continues to increase and the C/A ratio continues to decline. Zhou et al. [10] showed that when CaO%/SiO2% = 1.0–1.2, the Al2O3 content in the slag is greater than 6%, and compounds with high melting point are precipitated, which increases the melting point of the slag. At the same time, the titanium oxide (TiO2) produced by the flux–steel reaction easily reacts with (CaO) in the slag to form this high-melting phase (reaction (6)). This phase is likely to be CaTiO3 (perovskite), which precipitates at high temperatures [11]. Therefore, in the whole process of flux–steel reaction, the slag composition always changes in the direction that is conducive to the formation of high melting point inclusions at the slag interface. To verify the existence of high melting point CaTiO3 phase in thermodynamic calculations, XRD analysis was performed on the slag sample after the reaction, and the results are shown in Figure 7. From the XRD analysis results, it can be seen that the main phase compositions of the slag sample during the reaction process are CaTiO3, Fe elemental, and Ca2Al2SiO4, indicating that SiO2 and Fe2O3 in the slag sample have reacted with TiN or [Ti] in the steel to form Fe and TiO2 [12], and TiO2 combines with CaO to form high melting point perovskite. Therefore, the precipitation of high melting point inclusions leads to the increase of melting point and viscosity of slag.

(6) TiO 2 ( s ) + ( CaO ) = CaTiO 3 ( s ) ΔG ( 6 ) 0 = − 78636 .33 − 10 .80 T (J ⋅ mol − 1 ) .

Figure 6 Phase diagram of CaO–SiO2–Al2O3–Fe2O3–TiO2–Na2O (ωTiO2 = 1.23 mass%) (ωNa2O = 10.42 mass%) (ωFe2O3 = 1.13 mass%) obtained using FactSage 7.2.
Figure 6

Phase diagram of CaO–SiO2–Al2O3–Fe2O3–TiO2–Na2O (ωTiO2 = 1.23 mass%) (ωNa2O = 10.42 mass%) (ωFe2O3 = 1.13 mass%) obtained using FactSage 7.2.

Figure 7 XRD result of slag sample during the slag–steel reaction.
Figure 7

XRD result of slag sample during the slag–steel reaction.

From the changes in the composition content of the liquid steel and mold flux, it is evident that as the reaction progresses, the rate of the liquid steel and molten slag reaction continuously decreases, This phenomenon is associated with changes in the composition and viscosity of the mold flux, as well as an increase in solid-phase particles within the slag [13]. To study the influence of the change of chemical composition of slag on its viscosity after reaction, the viscosity module in FactSage 7.2 thermodynamic software was used to calculate the viscosity according to the chemical composition of A0–A5 mold flux samples extracted at different reaction times, with the target calculation temperature of slag of 1,550℃. According to the calculated data, the relationship between the chemical composition and viscosity of the slag sample was drawn, as shown in Figure 8.

Figure 8 Change of viscosity of slag with different reaction time.
Figure 8

Change of viscosity of slag with different reaction time.

As shown in Figure 8, with the progress of the reaction, the viscosity of the slag at the slag interface increased rapidly from 0.108 to 0.155. This may be due to the rapid accumulation of (Al2O3) and (TiO2) in the slag. Therefore, in order to study the reasons for the increase in viscosity, SEM analysis was conducted on the flux–steel interface, as shown in Figure 9(a). From the microscopic morphology of the flux–steel interface, it can be seen that a large number of irregularly granular TiO2·Al2O3 inclusions are mainly distributed near the flux–steel interface. According to the aforementioned thermodynamic analysis, this is formed by the reaction product TiO2 and Al2O3 in the slag. As shown in Figure 9(b), CaTiO3 in an irregular jagged shape is gradually precipitated from the slag phase at the interface, which is a high-melting precipitate produced by the reaction product TiO2 reacting with (CaO) in the slag [14]. Relevant studies show that the precipitation of a small amount of solid phase in slag significantly improves the viscosity near the interface [15]. Therefore, with the progress of the reaction, the contents of (Al2O3) and (TiO2) gradually increase, the formed CaTiO3 and TiO2·Al2O3 accumulate continuously at the reaction interface, and the inclusion layer is gradually formed at the flux–steel interface. The content of (SiO2) in the slag is gradually reduced, and the reactivity of the slag is weakened, resulting in the rapid increase of the viscosity of the slag. The diffusion of reactants and products in slag (SiO2) to the reaction interface is hindered, and the kinetic conditions of reactants in slag (SiO2) are worsened. Therefore, since the titanium–aluminum complex inclusions and perovskite minerals produced by the reaction are continuously deposited at the interface, the diffusion of components from slag to the reaction interface is limited, and the viscosity of mold flux near the flux-steel interface increases after the reaction.

Figure 9 SEM analysis of slag–steel interface: (a) SEM morphology of TiO2-Al2O3 and (b) SEM analysis of CaTiO3 (1 represents the steel phase, 2 represents the CaTiO3, and 3 represents the slag phase).
Figure 9

SEM analysis of slag–steel interface: (a) SEM morphology of TiO2-Al2O3 and (b) SEM analysis of CaTiO3 (1 represents the steel phase, 2 represents the CaTiO3, and 3 represents the slag phase).

Therefore, it can be seen that the slag steel reaction leads to an increase in viscosity and a decrease in mass transfer rate, which easily makes the solid particles and inclusions precipitated in the slag, as well as the reactive bubbles generated at the interface, more stable at the flux–steel interface, thereby greatly increasing the probability of steel surface defect formation after the reaction.

4.3 Effect of flux–steel reaction on the change of interface tension

4.3.1 Evaluation of surface tension of steel

In this study, the surface tension of liquid steel is calculated according to the following equation proposed by Tanaka et al. [16]:

(7) σ Fe = 1 , 910 − 825 log ( 1 + 210 mass % O ) − 540 log ( 1 + 185 mass % S ) ( mJ ⋅ m − 2 ) ( 0 ≤ mass % O / 0.0160 + mass % S / 0.0300 ≤ 1 ) .

Generally, equation (7) demonstrates that steel’s surface tension is determined primarily by O and S element contents [17]. In this current experiment, the oxygen content in the steel was measured using infrared absorption method, and the obtained results were applied to the equation to calculate the surface tension of the liquid steel. It was assumed that the sulfur content in the liquid steel remained constant throughout the experiment. As it was not possible to measure the aluminum content in situ during the experiment, therefore, the content of aluminum is determined by the chemical composition analysis of steel and slag samples taken from the furnace at different time intervals and rapidly quenched in water.

4.3.2 Evaluation of surface tension of mold flux

Methods for surface tension measurement include the maximum bubble pressure method and the ring method. For multicomponent mold flux, the following empirical formula can be used to calculate the surface tension of the mold flux [18]:

(8) σ Slag = ∫ 1 n X i K i , ( i = 1 , … n ) ,

where X i represents the molar fraction (%) of component i in the slag, while K i is the coefficient for component i. The values of K for different oxides can be found in the study by Wang et al. [19].

4.3.3 Calculation and variation relationship of interface tension between the steel and mold flux

Research reported experimental results on the interfacial tension changes between liquid iron alloy and molten CaO–SiO2–Al2O3 mold flux during chemical reactions by Suzuki et al. [20]. The changes in interfacial tension were determined using X-ray transmission microscopy through the sessile drop method. A theoretical calculation model for interfacial tension among steel with slag is based on a large number of scholars’ experimental results [21]. Subsequently, the following equation was derived to determine the interfacial tension:

(9) σ Fe / Slag = σ Fe 2 + σ Slag 2 − 2 ⋅ σ Fe ⋅ σ Slag ⋅ cos   θ 2 α ,

where σ Fe represents the surface tension of the steel in N/m, while σ Slag corresponds to the surface tension of the slag in N·m−1. The contact angle θ 2α between the steel and the slag is experimentally measured. The measurement data of contact angle are shown in Table 5. For this study, the surface tensions of σ Fe and σ Slag were derived from the model explained in detail in the preceding section. As a result, the interfacial tension σ steel/slag can be determined using the equation. First, the experimental results obtained from the interfacial interactions between the steel and the slag were used to calculate the changes in interfacial tension using the model equation mentioned above. The final calculation results are shown in Table 5, and their changes are shown in Figure 10.

Table 5

Changes of contact angle and interfacial tension of flux–steel with reaction time

T·(min−1) 0 5 10 15 20 30
θ (°) 53 32 38 41 42 42
σ (mJ·m−2) 1,380 1,320 1,335 1,340 1,342 1,345
Figure 10 Typical variation change in contact angle (L) and interfacial tension (R) between slag and steel reaction over time.
Figure 10

Typical variation change in contact angle (L) and interfacial tension (R) between slag and steel reaction over time.

Figure 10 illustrates the typical variation of interfacial tension between enameled steel and molten silicate slag over time. As shown in the figure, it can be observed that throughout the entire reaction process, the interfacial tension exhibits an overall decreasing trend. For reaction times of 5, 10, 15, 20, and 30 min, the surface tension between the steel and slag decreases from an initial value of 1,380–1,320 mJ·m−2, eventually stabilizing around 1,345 mJ·m−2.

According to Figure 10, it can be observed that throughout the entire reaction process, the interfacial tension between the steel and mold flux exhibits a rapid initial decrease from its initial value, followed by a minimum point, and eventually gradually increases towards an equilibrium state. Based on these results, the interfacial tension during the slag–steel reaction process can be described in detail through two reactions obtained from equation (10) (as shown in equations (11) and (12)), which mainly undergo the following two stages (①–② and ②–③):

(10) 4Al in Fe + SiO 2 in Slag = Al 2 O 3 in Slag + 3Si in Fe .

Stages ①–②: the decomposition of SiO2 in the slag into Si and O, entering the steel melt, leading to a rapid decrease in the interfacial tension:

(11) SiO 2 in Slag → Si in Fe + 2O in Fe .

Stages ②–③: the oxidation of Al in the steel melt, resulting in the precipitation of Al2O3 into the slag, causing a gradual increase in the interfacial tension and reaching equilibrium:

(12) 2Al in Fe + 3O in Fe → Al 2 O 3in Slag .

In the initial stage of the reaction, when the viscosity of the slag is low, the decomposition and dissolution of silicon dioxide (SiO2) occur immediately upon contact with the liquid steel. The silicon atoms diffuse into the liquid steel, whereas oxygen atoms prefer to adhere to the interface on the side of the liquid steel, since oxygen is a highly active element at interfaces [22]. The accumulation of oxygen atoms at the interface on the side of the liquid steel reduces the interfacial tension (represented by stages ①–② in Figure 10). With the progress of the reaction, the viscosity of the slag increases, and although the rate of oxygen diffusing from the interface to the slag phase slows down, the amount of oxygen atoms adsorbed on the interface decreases, resulting in the interfacial tension gradually increasing and tending to the final equilibrium state (as shown by stages ②–③ in Figure 10). Therefore, based on the experimental findings and theoretical analysis, it can be inferred that the primary factor influencing the alteration of interfacial tension in flux–steel is the redox reaction occurring at the interface.

5 Conclusion

The effect of slag–steel interface reactions on the surface quality of the cast slab was investigated using high-temperature simulation experiments and thermodynamic analysis. The impact of flux–steel reactions on the viscosity and interfacial tension of the slag was studied, and the following conclusions were drawn:

  1. The basic trends of component variations in the slag due to slag–steel interface reactions were observed. The content of (SiO2) in the slag continuously decreased from 35 to 25.10%, while the content of (Al2O3) increased from 5 to 9.95%. Correspondingly, the [Al] element in the liquid steel significantly decreased from 0.75 to 0.1% due to the occurrence of slag–steel reactions, whereas the variation in [Ti] content was minimal, indicating that [Al] reduction of (SiO2) is the predominant reaction. With increasing reaction time, the reaction rate gradually became more moderate.

  2. As the reaction progressed, flux–steel reactions led to the precipitation of solid phases with high melting points. The viscosity of slag gradually increased. At 300 s of reaction time, the viscosity rapidly increased from 0.108 to 0.138 Pa·s and then slowly elevated to 0.155 Pa·s. During the reaction process, the interfacial tension rapidly decreased from its initial value of 1,380 mJ·m−2 to a minimum of 1,320 mJ·m−2, and then it gradually increased to an equilibrium state.

  3. The reduction in interfacial tensions and the increase in viscosity caused by flux–steel reactions facilitated the accumulation of bubbles at the steel–slag interface, leading to foaming of the slag. This highly gas-laden foamy slag was more prone to being entrained into the liquid steel, resulting in difficulties in gas removal from the steel and the formation of surface porosity defects on the cast slab.

Acknowledgments

The authors are grateful for financial support from the cooperation project of School of Metallurgical Engineering, Anhui University of Technology and Baowu Ma Steel Co., Ltd, Fourth Rolling Steel Mill, (RD1720039).

  1. Funding information: This research was supported by the cooperative project of School of Metallurgical Engineering, Anhui University of Technology and Baowu Ma Steel Co., LTD, Fourth Rolling Steel Mill, (RD1720039).

  2. Author contributions: Zijian Huang and Qixuan Rui: writing-original draft, writing-review and editing, methodology, and formal analysis. Yingjiang Li and Xiaoguang Hu: writing-original draft, visualization, and project administration.

  3. Conflict of interest: Authors state no conflict of interest.

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Received: 2023-10-17
Revised: 2024-02-18
Accepted: 2024-02-20
Published Online: 2024-03-19

© 2024 the author(s), published by De Gruyter

This work is licensed under the Creative Commons Attribution 4.0 International License.

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https://doi.org/10.1515/htmp-2022-0327
Keywords for this article
surface defects; interface reaction; interfacial tension
Creative Commons
BY 4.0

Articles in the same Issue

  1. Research Articles
  2. De-chlorination of poly(vinyl) chloride using Fe2O3 and the improvement of chlorine fixing ratio in FeCl2 by SiO2 addition
  3. Reductive behavior of nickel and iron metallization in magnesian siliceous nickel laterite ores under the action of sulfur-bearing natural gas
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